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Effect of crack length-to-width ratio on crack resistance of high Cr-ODS steels at high temperature for fuel cladding application

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Oxide dispersion strengthened (ODS) steels with high Cr-content are extensively investigated in Europe, Japan and United States by the nuclear materials community for application to both advanced fission reactors and fusion systems. In comparison to standard high Cr-steels, the expected operation temperature range can be extended to 650 C or more because of their improved creep resistance. However, their crack resistance behavior in the high temperature range was less investigated. The aim of the present paper is to provide some insight on their fracture behavior at high temperature and different crack configurations, in particular shallow crack. Crack resistance measurements were performed on a 12%Cr-ODS steel using compact tension specimens at 650 C considering both shallow and deep crack configurations. Finite element calculations were performed on a typical fuel cladding tube geometry to assess the performances in terms of crack resistance. It is found that the temperature gradient across the wall should be maintained low enough to avoid cracking. After irradiation in corrosive environment, the boundary conditions might be further affected limiting therefore the lifetime of ODS cladding.
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  Effect of crack length-to-width ratio on crack resistance of high Cr-ODSsteels at high temperature for fuel cladding application R. Chaouadi ⇑ , M. Ramesh, S. Gavrilov SCK   CEN, Boeretang 200, 2400 Mol, Belgium a r t i c l e i n f o  Article history: Available online 21 February 2013 a b s t r a c t Oxide dispersion strengthened (ODS) steels with high Cr-content are extensively investigated in Europe, Japan and United States by the nuclear materials community for application to both advanced fissionreactors and fusion systems. In comparison to standard high Cr-steels, the expected operation tempera-ture range can be extended to 650   C or more because of their improved creep resistance. However, theircrack resistance behavior in the high temperature range was less investigated.The aim of the present paper is to provide some insight on their fracture behavior at high temperatureand different crack configurations, in particular shallow crack. Crack resistance measurements wereperformed on a 12%Cr-ODS steel using compact tension specimens at 650   C considering both shallowand deep crack configurations. Finite element calculations were performed on a typical fuel cladding tubegeometry to assess the performances in terms of crack resistance. It is found that the temperaturegradient across the wall should be maintained low enough to avoid cracking. After irradiation in corrosiveenvironment, the boundary conditions might be further affected limiting therefore the lifetime of ODScladding.   2013 Elsevier B.V. All rights reserved. 1. Introduction Among the candidate structural materials for fusion and fissionadvanced nuclear systems, oxide dispersion strengthened (ODS)steels are receiving considerable attention in Europe, Japan andin the US [1–6]. In particular, high Cr-ODS steels offer a numberof properties that are attractive for application at high temperatureunder intense irradiation [7–10]. However, the advantages of suchsteels such as their high strength at high temperature, their highcreep resistance and their relatively better irradiation resistanceare counterbalanced by their low fracture toughness leading tohigh ductile-to-brittle transition temperature (DBTT) [6,11–16]and their low crack resistance at high temperature [17,18].Recently, significant improvement of the DBTT was reached byrefining the oxide dispersion and better control of impurities andthe fabrication process [19–22]. Actually, the low fracture tough-ness at high temperature is more critical as these steels aresupposed to operate in the high temperature range where otheravailable non-ODS steels cannot be used. As a matter of fact, fewpreliminary results were indicating significantly low toughness athigh temperature [17,18]. It should be mentioned that althoughthe specimen orientation is not always reported, most of thepublished data are taken with the crack propagation occurringperpendicular to the extrusion direction, namely L–T. As it willbe seen later, this is not the weakest orientation and it is expectedthat the mechanical properties will be significantly poor when thecrack is orientated parallel to the extrusion direction.Within the EURATOM FP7 GETMAT project, three ODS steelswere selected for investigation with 9%Cr, 12%Cr and 14%Cr con-tent. The fracture toughness of the 12%Cr- and 14%Cr-ODS steelsat high temperature was found significantly low than commercial9%Cr steels [23]. The 9%Cr-ODS Eurofer steel reported in [18] exhibited the same tendency.From an application point of view, the measured fracturetoughness of high Cr ODS steels exclude them to be used as struc-tural materials of nuclear systems. This means that the fracturetoughness of high Cr ODS steels is too low to be used in thick com-ponents. The question remains whether they can be used for thinwalled components such as fuel cladding [2,24–31]. However, rep-resentative cracks in fuel cladding are shallow cracks that can befound on the inner or outer surface. Indeed, during the fabricationprocess, such penny-shape cracks can appear and it is interestingto assess their impact on the integrity of the fuel cladding duringoperation.In the following, we will experimentally measure the fractureresistance of the 12%Cr-ODS steel at 650   C in two crack configura-tions, deep (according to prevailing standards) and shallow crack,respectively. The tests will be performed in two specimen orienta-tions, L–T and T–L. Then, finite element calculations will beperformed on a typical thin wall cladding tube to estimate theactual loading during normal and accidental conditions to assess 0022-3115/$ - see front matter    2013 Elsevier B.V. All rights reserved.http://dx.doi.org/10.1016/j.jnucmat.2013.02.017 ⇑ Corresponding author. E-mail address:  rachid.chaouadi@sckcen.be (R. Chaouadi). Journal of Nuclear Materials 442 (2013) 425–433 Contents lists available at SciVerse ScienceDirect  Journal of Nuclear Materials journal homepage: www.elsevier.com/locate/jnucmat  the integrity of the fuel cladding using the measured fracturetoughness data. 2. Materials and experimental Within the GETMAT project, several 12%Cr-ODS steel plates(605  33  6.5 mm) fabricated by KOBELCO Research Institute in Japan were distributed among the participants for multi-disciplin-ary investigations. The chemical composition is given in Table 1.The final heat treatment consisted of annealing at 1150   C/1 h.The plate used is this work is labeled K11. Because of the sizelimitations of the available plate, the small compact tension (CT)geometry was selected. CT specimens were taken in two orienta-tions, L–T and T–L. The specimen size corresponds to  1 =  4  in.-CTspecimens with a thickness  B  = 6.25 mm.The  1 =  4  in.-CT specimens were precracked at room temperatureup to a crack length-to-width ratio of about 0.5 and further 20%-side grooved to ensure a uniform crack front propagation and topromote plane strain conditions. The stress intensity factor duringfatigue precracking was kept below about 18 MPa p  m at any time.For the deep crack configuration, the specimens in their stan-dard configuration, for example E1820 [32], were used. For theshallow crack configuration, the loading holes were located suchas  a / W   is between 0.1 and 0.2 (see Fig. 1). Depending on the usedholes for loading, either deep or shallow crack configuration can beexperienced. For illustration, in Fig. 2, the loading corresponds tothe shallow crack configuration. Tests were performed at 650   Cunder quasi-static loading rate of about 1 kJ/m 2 s and the  J  R  curveis determined using the energy normalization technique [33,34].At this high temperature, crack length monitoring with a clip gageand using the unloading compliance method cannot be used. Thedetails of the procedure can be found in [33,34]. The reasons of selecting the energy normalization method rather than a standard-ized method such as the normalization data reduction (NDR)technique [32] is its better accuracy and independence on experi-mental bias. Indeed, the method relies essentially on the specimendimensions, the load–displacement record and the measuredinitial and final crack lengths. An example is given in Fig. 3 wherethe energy normalization method is compared to three standardsmethods, namely the unloading compliance, the potential dropand the normalization data reduction.It should be noted that at 650   C, the tensile properties of the12%Cr-ODS steel exhibit a nearly perfectly plastic material withsignificant ductility anisotropy (see Fig. 4). As a matter of fact, thisanisotropy was reported by several authors including both tensileas well as fracture properties [35–40]. All latter authors reported atypical extruded microstructure with the grain elongated along theextrusion direction. 3. Testing and analysis For the deep crack configuration, the tests were duplicated tocheck material homogeneity and no significant difference wasfound [23]. Here, only a single test is reported as the aim is toassess the shallow versus deep configuration. The specimendimensions and testing conditions are given in Table 2.The  J  -integral is calculated according the ASTM E1820formulation [32]:  J   ¼  J  el  þ  J  pl  ¼  K  2  ð 1  V  2 Þ E   þ  g   U  pl B n ð W    a 0 Þ where  K   is the stress intensity factor,  E   is the Young’s modulus,  v  is the Poisson’s ratio,  W   is the specimen width,  B n  is the specimennet thickness,  a 0  is the initial crack length and  U  pl  is the areaunder the load–plastic displacement curve.Moreover,  g ¼ 2 : 0 þ 0 : 522 ð 1  a 0 = W  Þ . For crack resistance curve, the  J  -integral values are corrected forcrack growth according to:  J  ð i Þ  ¼ K  2 ð i Þ   ð 1  V  2 Þ E  þ  J  pl ð i  1 Þ  þ g ð i  1 Þ ð W    a ð i  1 Þ Þ U  pl ð i Þ   U  pl ð i  1 Þ B n    1   c ð i  1 Þ a ð i Þ   a ð i  1 Þ W    a ð i  1 Þ   where  g ð i  1 Þ  ¼ 2 : 0 þ 0 : 522 ð 1  a ð i  1 Þ = W  Þ  and  c ð i  1 Þ  ¼ 1 : 0 þ 0 : 76 ð 1  a ð i  1 Þ = W  Þ . However, the  g  and  c  factors given above are only valid for deepcrack configuration, namely, CT specimens with a crack-to-widthratio, a / W  , close to 0.5. In the case of shallow cracks, these formulasmight not be valid. Therefore, 3D finite element calculations wereconducted by Ramesh [41] on a 1/2T-CT specimen with a crack-to-with ratio varying between 0.1 and 0.8 to derive the  g -factor. Notethat for crack extension levels under consideration, the crackgrowth correction accounted by the  c -factor remains close to 1.Therefore, only the  g -factor was determined as a function of cracklength-to-width ratio.It is important to mention that, for these finite element simula-tions, the compact tension geometry was adapted to avoid largedeformation at the pin holes. Therefore, for shallow cracks, typi-cally  a / W   < 0.35, the geometry was modified by increasing thematerial thickness around the loading pin holes and consequentlyavoiding excessive deformation of the pin holes. Note that modifi-cation holds only for the finite element simulation because atougher material was used in those simulations. In the case of ODS steels, the crack resistance is so low that no excessive defor-mation occurred at the loading pin holes.The finite element results can be summarized as: g  ¼  2 : 72  0 : 52 ð a = W  Þ  3 : 75 ð 1  a = W  Þ 6 It should be mentioned that by moving the pin hole position to re-duce the crack length-to-width ratio, the loading changes from pre-dominantly bending for deep crack to predominantly tensile forshortcrack. Therefore,itis interestingto comparethe finiteelementresults obtained by Ramesh [41] to results obtained on single edgecrack tension, SE(T), specimens obtained by [42,43]. The compari-son is shown in Fig. 5 whichindicates a consistent trend, supportingthe use of the above derived equation for shallow crack configura-tion while the ASTM formula is applied for the deep crackconfiguration. 4. Test results Four 1 =  4  in.-CT specimenswere tested at 650   C in two crackcon-figurations (shallow crack versus deep crack) in two orientations(L–T versus T–L). The specimen dimensions and testing conditionsare given in Table 2.As it can be seen, a crack extension of 1 mm was reached byapplying a load equivalent to a  J  -integral level of only 3.5 kJ/m 2 for the deep crack CT in the T–L orientation. This is significantlylow and lower than the value in the L–T orientation in which55 kJ/m 2 were required to extend the crack by 0.77 mm. Similarly,for the shallow crack configuration, almost 3 mm crack extensionwere reached with an equivalent load of about 35 kJ/m 2 in the  Table 1 Chemical composition (in wt%). Cr W Ti Si Ni Mn Y  2 O 3 11.59 1.87 0.22 0.10 <0.01 <0.01 0.23426  R. Chaouadi et al./Journal of Nuclear Materials 442 (2013) 425–433  T–L orientation while only 1 mm was achieved with a load corre-sponding to 184 kJ/m 2 in the L–T orientation.For each specimen, the crack resistance curve is determinedfrom which the engineering crack initiation toughness establishedat 0.2 mm crack extension,  J  0.2mm , the tearing modulus defined as T   ¼  dJ da   E  ð 1  v  2 Þ r 2  y and the elastic–plastic stress intensity factor K   J  c   ¼  ffiffiffiffiffiffiffiffiffiffiffiffi   J  0 : 2mm  E  ð 1  v  2 Þ q   are derived. The results are shown in Table 3 for allspecimens. As expected, the stress intensity factor  K   J  c   of the deepcrack configuration is only 21 MPa p  m in the T–L orientation. Thisis the reason why such a material cannot be used as a structuralmaterial. For the shallow crack configuration in the T–L orienta-tion, the stress intensity factor is enhanced to 46 MPa p  m. Thecrack resistance curves,  J  - D a , of the various specimens are showninFig.6whilethecharacteristicparametersaredepictedinTable3. As it can be seen, this ODS steel exhibits a significant crackresistance anisotropy and in particular, a significantly weak crackresistance in the extrusion direction. In terms of stress intensityfactor, the results are shown in Fig. 7. The low fracture resistancein the T–L orientation is very critical as longitudinal cracks canbe generated during the extrusion process. Before investigatingthe consequences of such low crack resistance in a component, itis interesting to examine the cracking behavior and crack surfaceusing microscopic techniques.One of the characteristics of the 12%Cr-ODS steel cracking canbe evidenced by optical microscope observation of the crack Fig. 1.  CT-specimen configuration allowing deep ( a / W   = 0.5) and shallow crack ( a / W   = 0.1) configurations. Fig. 2.  Shallow crack CT-specimen testing. Fig. 3.  Typical example of crack resistance curve determination proceduresaccording to various methods as applied to a 508 Cl.3 steel: unloading compliance,potential drop, normalization data reduction and energy normalization techniques. Fig. 4.  Tensile properties of the 12%Cr-ODS steel at 650   C in the L- and T-orientation [23]. R. Chaouadi et al./Journal of Nuclear Materials 442 (2013) 425–433  427  propagation ahead of the fatigue precrack. The energy required toinitiate and propagate the crack is so low that the crack extensionoccurs with very small crack tip opening displacement (see Fig. 8).Actually, the final crack propagation (Fig. 8c) is very similar to a fa-tigue precrack (Fig. 8a) for which the applied stress intensity factordoes not exceed 18 MPa p  m. Similar observations were reported byAlinger et al. [35]. Moreover, the crack propagation occurs on thesame plane as the fatigue precrack without deviation or zigzaggingas usually observed on more ductile materials. This clearly indi-cates the very low crack resistance of this ODS steel.The effect of specimen orientation on the crack resistance wasalso examined using scanning electron microscopy (SEM) observa-tions. For L–T specimens, we can notice secondary cracks appear-ing on the fracture surface but that are perpendicularly orientedwith respect to the crack plane (see Fig. 9a). Actually, while the testis such as to favor crack extension in the T-direction, theses crackspropagate in the L-direction (extrusion direction). On the T–L spec-imens in which crack propagates in the L-direction, such secondarycracks are not observed (see Fig. 9b). This clearly demonstrates thelow crack resistance of the 12%Cr-ODS steel when the crack ispropagating in the extrusion direction.Large SEM magnification of the fracture surface in the region of interest shows that in the weakest orientation, namely T–L, thegrain structure is elongated along the extrusion direction (seeFig. 10). Although not unambiguously observed, it is likely thatthe fracture is intergranular resulting from grain boundary de-cohesion. We performed stereo SEM micrographic observationsthat show that while in the L–T orientation, the fracture surfaceexhibits a typical dimple structure with a tortuous fracturelandscape, the T–L orientation shows a much more flatter surfacewith textured structure and intergranular aspect (see Fig. 10). Notethat in the high temperature regime, intergranular cracking wasalso reported by some authors [44–47].The small oxide particles and other second phase particles(inclusions) are definitely playing an important role in offeringpreferential sites for void nucleation at the interface particle–matrix. The deformation incompatibility between the hard oxideparticles and the soft matrix favors easy void nucleation and theircoalescence is facilitated by the high density of oxide particles. As aresult, very limited energy is required to initiate and further prop-agate a crack. Due to the extrusion process, it is possible that thesesecond phase particles (oxides, inclusions) are aligned in coloniesoriented in the extrusion direction that significantly decrease thecrack resistance. According to Alinger [35], the low toughnesswhen the crack is parallel to the L-direction is due to the presenceof a large volume fraction of alumina stringers resulting fromimpurity contamination. Further microstructural investigationswill be required to better identify all the features that areresponsible of the low crack resistance and anisotropy of the ODSsteel. In any case, the fractureproperties of ODS steels, in particularin the weakest direction, need to be improved to a level that do not jeopardize their usability. 5. Finite element simulation of cracked thin wall cladding tube One of the potential application of high Cr-ODS steels is for thefuel cladding. For high burnup fuel cladding, ODS steels are consid-ered as most promising [30].In Japan, a number of authors [27,30,47] reported successfulfabrication of high Cr ODS cladding tubes. The creep rupturestrength of a 9%Cr-ODS steel with 0.3%Y  2 O 3  at 700   C was signifi-cantly increased. A number of fuel pins made of high Cr-ODS areunder irradiation in the BOR-60 reactor [30,31] at high tempera-ture, 650 and 700   C with promising results, i.e., no fuel pin failureobserved, were obtained after irradiation to 21 DPa. Results athigher neutron dose are not reported yet.  Table 2 Specimen dimensions and testing conditions of the 12%Cr-ODS steel at 650   C. Specimen Orientation Width(mm)Thickness(mm)Net thickness(mm)Crack length(mm)Crack length to widthratioCrack extension(mm)  J   at end of test (kJ/m 2 ) W B B n  a 0  a 0 / W   D a end  J  end K11-32 T–L 7.66 6.17 4.86 1.08 0.14 2.91 33.0K11-27 L–T 7.52 6.22 4.87 1.00 0.13 0.98 184.2K11-24 T–L 12.51 6.25 4.97 6.66 0.53 1.05 3.5K11-20 L–T 12.50 6.26 5.12 6.61 0.53 0.77 54.8 Fig. 5.  Effect of crack configuration (shallow to deep) on the plastic  g  factor of specimens submitted to tensile, bending and combined tensile/bending loading. Allsymbols are from finite element calculations.  Table 3 Crack resistance test results of the 12%Cr-ODS steel at 650   C. Specimen Orientation Crack length towidth ratio (–)  J   at onset of crackextension (kJ/m 2 )Tearing resistance(kJ/m 2 p  mm)  J   at 0.2 mm crackextension (kJ/m 2 )Tearingmodulus (–)Stress intensity factor(MPa p  m) a 0/ W J  i  J  t   J  0.2mm  T  K11-32 T–L 0.14 3.6 17.3 11.3 32 46K11-27 L–T 0.13 13.1 172.5 90.2 315 129K11-24 T–L 0.53 1.5 2.0 2.35 3.7 21K11-20 L–T 0.53 9.3 51.9 32.5 94.9 77428  R. Chaouadi et al./Journal of Nuclear Materials 442 (2013) 425–433
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